ce -/ 7 ^ j - z ^7
NASA Contractor Report 172237
^9840003226
MATERIAL CHARACTERIZATION OF STRUCTURAL
ADHESIVES IN THE LAP SHEAR MODE
Steven C. Schenck and Erol Sancaktar
CLARKSON COLLEGE OF TECHNOLOGY
Potsdam, New York 13676
Grant NAGl-284 ^ c i .. . - .
September 1983
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National Aeronautics and
Space Adnninistration
Langley Research Center
Hampton, Virginia 23665
ACKNOWLEDGMENTS
This project was sponsored by NASA's Langley Research Center under
the NASA Grant NAG- 1-284. Dr. William S. Johnson was the NASA
6
technical monitor and his assistance is gratefully acknowledged.
Thanks are also due to Dr. Terry L. St. Clair of NASA Langley for his
friendly guidance and assistance throughout the implementation of this
project.
This work constituted Steven C. Schenck's M.S. Thesis.
ii
NOMENCLATURE
A,B,C
A' ,B',C
A'',B",C'
} Cfp
G
H(t)
K,n
R
T
YtY
E V
iJ ij
0
U
T,T
'^ult
T
O
T(t)
X
Y
o
material constants for Maxwell model
material constants ^
material constants for Kelvin model
functions of temperature
shear elastic modulus
unit step function
material constants
universal gas constant
time constant
rupture time
absolute temperature
constant activation energy per mole
one dimensional shear strain and shear strain rate
elastic and viscoelastic strains
elastic limit stress
coefficient of viscosity in shear
one dimensional shear stress and shear stress rate
ultimate shear stress
level of constant stress
material constants
time dependent shear stress
elastic limit shear strain
time dependent material property
level of initial (instantaneous) strain in a creep test
V
Poisson’s ratio
LIST OF FIGURES
Figure
Page
1
Clip-on Gage Attachment to the Single Lap Specimen
for Adhesive Deformation Measurement.
19
2
Constant Strain Rate Stress-Strain Behavior of
FM 73 and Comparison with Theory.
22
3
Variation of Ultimate Shear Stress with Initial
Elastic Strain Rate and Comparison with Ludwik's
Equation for FM 73 Adhesive.
24
4
Variation of Maximum Shear Strain with Initial
Elastic Strain Rate for FM 73 Adhesive.
26
5
Variation of Elastic Limit Shear Stress with Initial
Elastic Strain Rate and Comparison with Ludwik's
Equation for FM 73 Adhesive.
27
6
The Effects of Temperature on the Constant Strain
Rate Stress-Strain Behavior of FM 73 Adhesive.
28
7
Room Temperature Creep Response of FM 73 Adhesive.
29
8
Creep Behavior of FM 73 Adhesive at 130°F(54°C)
Condition.
30
9
Creep Behavior of FM 73 Adhesive at 180°F(82°C)
Condition.
31
10
Creep Rupture Data and Comparison with Zhurkov's
Equation for FM 73 Adhesive.
33
11
Creep Rupture Data and Comparison with Crochet's
Equation for FM 73 Adhesive.
34
12
Variation of Maximm (Safe) Creep Stress with
Environmental Temperature for FM 73 Adhesive.
36
13
Constant Strain Rate Stress-Strain Behavior of
FM 300 Adhesive and Comparison with Theory.
37
14
Variation of Ultimate Shear Stress with Initial
Elastic Strain Rate and Comparison with Ludwik's
Equation for FM 300 Adhesive.
38
15
Variation of Maximum Shear Strain with initial
Elastic Strain Rate for FM 300 Adhesive.
40
Figure Page
16 Variation of Elastic Limit Shear Stress with Initial 41
Elastic Strain Rate and Comparison with Ludwik's
Equation for FM 300 Adhesive.
17 The Effects of Temperature on the Constant Strain 42
Rate Stress-Strain Behavior of FM 300 Adhesive.
18 Room Temperature Creep Response of FM 300 Adhesive. 43
19 Creep Behavior of FM 300 Adhesive at 180°F(82®C) 44
Condition.
20 Creep Behavior of FM 300 Adhesive at 250°F(12l°C) 45
Condition.
21 Creep Behavior of FM 300 Adhesive at 30 CPf(149°C) 46
Condition.
22 Creep Rupture Data and Comparison with Zhurkov's 48
Equation for FM 300 Adhesive.
23 Creep Rupture Data and Comparison with Crochet's 49
Equation for FM 300 Adhesive.
24 Variation of Maximum (Safe) Creep Stress with 50
Environmental Temperature for FM 300 Adhesive.
25 Constant Strain Rate Stress-Strain Behavior of 51
Thermoplastic Polyimidesulfone Adhesive and
Comparison with Theory.
26 Variation of Ultimate Shear Stress with Initial 53
Elastic Strain Rate and Comparison with Ludwik's
Equation for Thermoplastic Polyimidesulfone Adhesive.
27 Variation of Maxirntm Shear Strain with Initial 54
Elastic Strain Rate for Thermoplastic
Polyimidesulfone Adhesive.
28 The Effects of Temperature on the Constant Strain 55
Rate Stress-Strain Behavior of Thermoplastic
Polyimidesulfone Adhesive.
29 Room Temperature Creep Response of Thermoplastic 56
Polyimidesulfone Adhesive.
30 Creep Behavior of Thermoplastic Polyimidesulfone 58
Adhesive at 250°F(121°C) Condition.
V
Page
Figure
31 Creep Behavior of Thermoplastic Folyimidesulfone 59
Adhesive at 350°F(177°C) Condition.
32 Creep Behavior of Thermoplastic Folyimidesulfone 60
Adhesive at 450°F(232°C) Condition.
33 Typical Fracture Surfaces of Titanium-Polyimidesulfone 61
Specimens Exhibiting the Effects of Increasing
Temperature.
34 Creep Rupture Data and Comparison with Zhurkov's 62
Equation for Thermoplastic Folyimidesulfone Adhesive.
35 Creep Rupture Data and Comparison with Crochet's 64
Equation for Thermoplastic Folyimidesulfone Adhesive.
36 Variation of Maximum (Safe) Creep Stress with 65
Environmental Temperature for Thermoplastic
Folyimidesulfone Adhesive.
vi
LIST OF TABLES
Table
Page
1
Mechanical Properties of FM 73, FM 300 and
Thermoplastic Polyimidesulfone Adhesives at
Comparable Testing Conditions.
66
B-2
Dimensions for Single Lap Specimens Bonded vith
FM 73 Adhesive.
78
B-3
Dimensions for Single Lap Specimens Bonded with
FM 300 Adhesive.
79
B-4
Dimensions for Single Lap Specimens Bonded with
Thermoplastic Polyimidesulfone Adhesive.
80
vii
CHAPTER 1
INTRODUCTION
Structural adhesives are preferred over the customary penetration
techniques for joining of modern lightweight-composite materials.
Their usage increased considerably during the last decade in the
aerospace, automotive, naval, and many other industries. Structural
adhesives are superior to conventional means of joining materials
because of their flexibility and toughness, lightness of weight,
exceptional thermal stability, solvent and moisture resistance, and
excellent mechanical properties at room and elevated temperatures.
In joining mechanical components, the structural adhesive and
mode of bonding (i.e. lap, butt, scarf, etc.) to be used is usually
determined by in— service requirements. Single lap joints are widely
used due to their applicability in many industrial designs.
In the absence of catastrophic crack propagations, the failure of
lap joints may occur in one of the following modes.
1) Rupture of the adhesive layer, when the ultimate stress is
reached;
2) Creep rupture of the adhesive layer when a high level of
constant load is used;
3) Failure of the adherends.
Failure in the third mode will be unlikely when metal adherends such
as steel or titanium are used. It is necessary, however, to consider
the first two modes of failure for the design of adhesively bonded
joints.
For design purposes, one often needs only the elastic properties
1
(namely Young's modulus and elastic limit stress) and failure stresses
as a function of rate, time and temperature. If such is the case,
then a satisfactory characterization can be obtained with the use of a
semi-empirical approach to describe the failure stresses as a function
of rate and temperature for constant strain rate loading and as a
function of time and temperature for constant load conditions. This
approach assumes that the viscoelastic effects are negligible in the
initial portion of the stress-strain curve, so that an initial elastic
strain rate can be defined. Previous work on a variety of polymeric
materials and adhesives proved this assxjmption to be a valid one [1,
2 , 3 , 4 , 5 ].
If information on the magnitude of strains (passed the elastic
limit) is also required, then one needs to use a theoretical approach
with the application of a mechanical model to characterize the
material. Information on failure stresses can be extracted from such
an approach if perfectly plastic flow is present (i.e. if the constant
strain rate stress-strain curve has an asymptote). However, one still
needs to use semi-empirical methods associated with the model, to
obtain the creep rupture stresses. Furthermore, the use of a
semi-empirical method is necessary to obtain rupture stresses for
constant strain rate conditions, when perfectly plastic flow is not
present.
It should also be noted that for some linear thermoplastic
adhesives the effects of rate on the failure stresses may be
negligible enough to permit the use of a nonlinear elastic relation to
characterize the constant strain rate stress-strain behavior.
The semi-empirical relation proposed to describe rupture stresses
2
at room and elevated temperatures are Ludwik's equation [6] for the
constant strain rate condition and Zhurkov's [7] and Crochet's [8]
equations for the creep condition. It should be noted that the
original form of Ludwik's and Crochet's equations do not include
thermal effects. This report proposes an empirical modification of
Ludwik's equation and an empirical extension of Crochet's equation to
describe the effects of temperature*
Literature Review
The strength of adhesively bonded joints are known to depend on
rate, time and temperature* H. F* Brinson [1] demonstrated that
yielding in polycarbonate is both rate and time dependent* H. F*
Brinson et* al* [2] studied the stress-strain and rate and time
dependent yield behavior of Metlbond adhesive in the bulk form.
Uniaxial tensile constant strain rate, creep and relaxation tests were
performed for this purpose* According to Brinson, the rate dependent
variation of the linear elastic limit and the failure stresses for the
bulk adhesive could be described with a semi-empirical equation
proposed by Ludwik [6] , whereas time dependent variation of yielding
could be correlated with a delayed yield equation proposed by Crochet
[8] . Furthermore, the modified Bingham model was shown to represent
adequately the constant strain rate stress-strain behavior of the bulk
adhesive* The same model was later used by Sancaktar [3] for
characterizing the viscoelastic shear behavior of a structural
adhesive in the bulk and bonded forms.
Sancaktar and Padgilwar [5] studied the rate and time dependent
mechanical behavior of LARC-3 adhesive in the lap shear mode by using
3
constant strain rate and creep tests. The constant strain rate
stress-strain curves of LARC-3 contained regions of elastic behavior
followed by regions of viscoplastic b^^havior. MaxiTnum and elastic
limit shear stresses were shown to be rate dependent and their
variations with the initial elastic strain rate agreed well with the
semi-empirical equation proposed by Ludwik [6] . A creep to failure
phenomenon was shown to exist and was correlated with the delayed
yield equation proposed by Crochet [8] . Analytical predictions based
on the Chase-Goldsmith model [4] were shown to agree well with the
experimental stress-strain-strain rate data.
Temperature also has a strong influence on the mechanical
stress-strain properties of adhesives. For instance, the effects of
temperature on the strength of lap shear specimens bonded with a high
molecular weight thermoplastic adhesive have been investigated by
Bugel, Norwalk and Snedeker [9]. They reported a sharp decrease in
the joint strength for temperatures above 160^F (71°C).
Some mechanical data on the Thermoplastic Polyimidesulfone
adhesive is available in a paper by St. Clair and Yamaki [10]. They
report that the lap shear strength decreased from 4150 psi (28.6 MPa)
to 2620 psi (18.1 MPa) when the test temperature was increased from
the room condition to 450^F (232^C). St. Clair and Yamaki also report
that thermal aging of the adhesive in the bonded form at temperatures
up to 450 ^F (232^C) resulted in a reduced lap shear strength. For
example, the ambient temperature strength of lap shear specimens that
had been aged for 5000 hours at 450^F (232^0 was 3640 psi (25.1 MPa),
while the same strength for the unaged specimens was 4150 psi (28.6
MPa).
4
Adhesive manufacturers generally do not supply detailed
information on the rate, time and temperature dependent variation of
the lap shear properties. When data is available, it is usually
limited to the ultimate strength valuers.
Experimentally measured lap joint strength values for
Thermoplastic Polyimidesulfone , FM 73 and FM 300 adhesives are
available in the literature [10, 11, 12]. Information in regard to
the rate, time and temperature dependent single lap mechanical
behavior for these three adhesives, however, is yet to be obtained
during this investigation.
The present investigation uses single lap specimens which were
supplied by the NASA Langley Research Center. For such specimens the
shear stresses are calculated as load divided by the overlap area, as
prescribed by the ASTM standards. It should be noted, however, that
the shear stresses calculated in this manner provide only approximate
values as the actual shear stress distribution along the overlap area
is not uniform but in fact is part of a biaxial stress state. The
exact form of the stress distribution in single lap joints has been
described by Goland and Reissner [l3] .
Objective
The purpose of the current investigation was to find a practical
method for characterizing structural adhesives in the bonded lap shear
mode. The applicability of the proposed characterization method was
to be assessed by:
1) obtaining experimental data on the delayed failure( creep to
yield) behavior of three different model adhesives in the bonded
5
form at both room and elevated temperatures;
2) obtaining experimental data on the constant strain rate shear
stress-strain behavior of three different model adhesives in the
bonded form at both room and elevated temperatures;
3) developing or identifying viscoelastic or nonlinear elastic
models which will describe the mechanical behavior observed in
the above mentioned experimental data;
4) determining whether Crochet's equation will describe the
delayed failure behavior observed in experiments;
5) assessing the applicability of Zhurkov's (modified) equation
in describing the time and temperature dependent behavior of
failure stresses under constant loading conditions;
6) determining whether Ludwik's equation will describe the rate
dependency of rupture stresses and elastic limit stresses and
strains.
The three model adhesives, one of which was developed at the NASA
Langley Research Center, were applied on titanium adherends in the
form of single lap joints to represent in-service conditions. The
adhesives are;
1) Thermoplastic Polyimidesulfone
2) FM 73, and
3) FM 300 adhesives.
1) Thermoplastic Polyimidesulfone is a novel adhesive developed at
NASA - Langley Research Center. It has thermoplastic properties
and solvent resistance. It is processable in the 482 - 662 F®(250
O
- 350 C) range and has high temperature resistance [lOJ .
6
2) FM 73 is a modified epoxy adhesive film with carrier cloth. It is
manufactured by the American Cyanamid Company. Its product
information brochure reports a service temperature range of ~67^F
to 248°F (-55°C to 120 °C) with high moisture resistance til].
3) FM 300 is a modified epoxy adhesive film supplied with a "tricot -
knit" carrier cloth that offers a good mixture of structural and
handling properties. It is manufactured by the American Cyanamid
Company as a supported film. Its product information brochure
reports a service range of -67^F to 300°F (-55^C to ISO^C) with
excellent moisture and corrosion resistance [l2l .
The second chapter of this report is devoted to the analytical
description of adhesive mechanical behavior with the presentation of
constitutive equations and their solutions for constant strain rate
and load conditions.
Chapter 3 outlines the experimental procedure. Materials
selected for the study and specimen features are reviewed.
Chapter 4 is devoted to the discussion of the results.
Theoretical results obtained from the proposed constitutive equations
are compared with the experimental data.
The final chapter presents the conclusions.
7
CHAPTER 2
ANALYTICAL CONSIDERATIONS
In designing single lap joints bonded with a viscoelastic
adhesive, a complete material characterization of the adhesive is
needed. Such a characterization usually includes the elastic
properties (namely Young's Modulus and elastic limit stress) and the
failure stresses as a function of strain rate, time and temperature,
A satisfactory characterization can be obtained by describing the
rupture stresses as a function of rate under constant strain rate
loading and as a function of time and temperature under constant
loading by using a semi-empirical approach. In this approach, the
rate dependency of the rupture stresses under constant strain rate
loading can be expressed with the semi— empirical equation proposed for
metals by Ludwik (as reported by Thorkildsen [6]) in the form
e
T “ t' + t" Log (J— ) ,
ult
o
where t - is the ultimate shear stress, Y is the initial elastic
ult
o
shear strain rate and x*, x*’ and Y are material constants* This
equation was used successfully by Brinson et* al* [1,2,3] and
Sancaktar [5] in describing the rate dependence of rupture stresses
for many polymeric materials in the bulk tensile shear, and bonded lap
shear modes.
The same form of equation (1) can also be used to describe the
variation of elastic limit shear stresses and strains with initial
elastic shear strain rates. These expressions may be written as
8
0 = e’ + e” Log(Y/Y')
( 2 )
(j) = <J)' + (J)'' Log(Y/v')
where additional material constants are defined accordingly.
Superposition of temperature effects on equations (1), (2) and
(3) would require experimental justification. For example,
examination of rate vs. strength data for copper [15] at various
temperatures (73°F < T < 1832®F, 23®C < T < 1000°C) indicates a near
parallel shift in the strain rate vs. strength relation particularly
in the 392°F (200®C) to 1472®F (800 ®C) region. Such a parallel shift
in equations (1), (2) and (3) could be expressed as
•Tuit = {a^}[x’]+ t" Log (y/y') (4)
0 = {b^}[0l + 0" Log (y/y') ^5)
(j) = ']+<!>" Log (y/y') (6)
where a,p, brp and c,p are functions of temperature.
Temperature dependent delayed failure of structural adhesives can
be described with the use of Zhurkov's and/or Crochet's creep-rupture
equations. The modified (by Slonimskii [16D form of Zhurkov's
creep-rupture equation is given by:
tr - to exp [^ - YCt] O)
9
where
= rupture time
tQ = a constant normally about lo” (stated to be the period
of natural oscillation) “
Uq = a constant activation energy per mole (associated
as an energy-barrier term)
Y = a constant
o' r applied uniaxial stress
R r universal gas constant
T r absolute temperature.
This equation was recently used by Brinson et. al. to describe the
creep-rupture behavior of various polymer-matrix composites [171.
A decaying exponential relation of the form
o(t) “A' + B' exp (-C'x) (8)
was proposed by Crochet to describe delayed failures. In equation
(8), oKt) is the time dependent maximum stress. A', B' and C are
material constants and X is a time dependent material property given
by
r / V E . > V E . ,Js
(9)
V E
where and refer to viscoelastic and elastic strains,
respectively. Equation (8) can be interpreted for the pure shear
condition. Use of a viscoelastic model is necessary, however, in
order to obtain the material property X. For practical purposes.
10
simple linear viscoelastic models, Maxwell fluid or Kelvin solid can
be used to facilitate the solution to equation (9). The use of
Maxwell model results in
s
Ct
. T(t) « A + B exp [- -jp t] (10)
and the use of Kelvin model results in
T(t) - A" + B” exp [C"e ^
where :
x(t) “ time dependent maximum shear (rupture) stress
T = level of constant stress,
o
VI = coefficient ,5>f viscosity in shear
G “ modulus of elasticity in shear
Apparently equation (11) describes a much faster approach in time
of the rupture stresses to their asymptotic values.
The asymptotic values A and (A”+B") of equations (10) and (11)
respectively, are important design parameters, as they represent the
maximum safe stress values below which creep failures are not expected
to occur. One can express A and/or (A"+B”) values as a function of
the applied temperature to represent experimental data in an empirical
fashion.
In order to predict the stress- strain behavior of the adhesive
and to extract failure information for constant strain rate and
constant stress modes, one needs to determine the viscoelastic model
which will describe the material adequately. If the rate effects on
11
the material are determined to be negligible, then the use of a
nonlinear elastic relation may also be adequate. Brinson et. al. lU
used the modified Bingham model (shown in Figure 2), which was
modified from the Schwedoff model [18] for characterizing the tensile
behavior of Metlbond adhesive. The modified Bingham model was later
used by Sancaktar [3] for characterizing the shear behavior of the
same adhesive. For that work, the modified Bingham model was applied
in the form,
T " G Y ^ ®
Y
( 12 )
where t and Y are shear stress and strain respectively, and 6 is the
elastic limit shear stress*
To obtain the constant strain rate relation, the above
constitutive equation is solved according to the condition
o
. Y “ R “ constant ( 13 )
to result in
• T - G Y X <_ 0 • (14)
T - 0 + y Y [1 - e*P ® ^ ^ -"^ult (15)
where ^ is the elastic limit strain.
The creep relation is obtained by using a step loading
12
(16)
T(t) = H(t)
where H(t) represents the unit step function and is the level of
constant stress, with the initial condition
Y(t-o) -
to obtain
Y(t)
T -0
o
y
(18)
In order to extract creep-rupture information from the given
model, Crochet's delayed failure equation (equation 8) is used.
Equations (8) and (9) can be applied to the shear condition by
subtracting the elastic strain
T /G
o o
(19)
from the model's creep equation (equation 18) , to obtain x
substituting into equation (8) to result in
i(t) « A + B exp [-K(T(t)-0)t] .
( 20 )
Equation (20) is the creep-rupture equation based on the modified
Bingham model. It should be noted that the time— temperature
superposition principle can be used along with equation (20) to make
13
it useful at different temperature levels. Naturally, it is not
expected that the modified Bingham model will characterize all
available adhesives. One can, however, use other viscoelastic models
and follow the procedure described above to characterize most
adhesives. Some examples of such application would be the use of the
Chase-Goldsmith model by Sancaktar [5] to characterize LARC-3 adhesive
in the bonded form and the use of Schapery's nonlinear constitutive
equation [19] by Brinson et. al. [20] to characterize SMC-25 fiber
reinforced polyester composite.
If the rate effects on adhesive behavior is determined to be
weak, then a nonlinear elastic relation can be used to fit the
stress-strain data. The power function relation expressed for the
state of pure shear in the form:
T - K (21)
where K and n are material constants, is proposed for this purpose.
14
CHAPTER 3
EXPERIMENTAL PROCEDURES
Materials and Specimen Features ^
The processing of the test specimens bonded with Thermoplastic
Polyimidesulfone has been described in detail by St. Clair [lO]. Two
one-inch (2.54 cm) wide. 0.05 inch (0.127 cm) thick titanium alloy
strips which are exact copies of each other are sand blasted with
aluminum oxide of 120 mesh. These duplicate strips are then coated
with poly (amide-acid sulfone) which is formed by mixing 0.0569 lbs
(25.8 gms) of 3,3' ,4,4'-Benzophenonetetracarboxylic dianhydride(BTDA)
in a solution made up of 0.0439 lbs (19.9 gms) of
3.3'-diaminodiphenylsulfone(3,3'-DDS) in 0.5701 lbs (258.6 gms) of bis
(2-methoxyethyl) ether. Fifteen minutes of thermal treatments at 212
°F (100*^0 and 392°F (200 °C), respectively, are applied to convert the
amide-acid to the amide by removing the solvent. The adhesive layers
are then applied onto the adherend strips. A 0.004 inch (0.010 cm)
thick piece of woven glass cloth is inserted between the strips to
control the adhesive thickness. A pressure of 200 psi (1.38 MPa) is
applied and the specimen heated at a rate of 9°F/min (5°C/min) up to
572°F (300 °C). The bonded specimen is then cooled and removed.
The adhesive, FM 73, is a modified epoxy adhesive film with a
polyester knit fabric that offers optimum physical properties. It is
manufactured by the American Cyanamid Company. Its product
information brochure reports a service range of -67 ‘V to 248 F (-55 C
to 120°C) with a high resistance to moisture.
The bonding procedure with FM 73 adhesive is described in detail
15
by its product information brochure [11] • Two duplicate one~inch
(2.54 cm) wide, 0.05 inch (0.127 cm) thick titanium alloy strips are
cleaned and dried to provide a grease-free surface for bonding.
Patterns of FM 73 adhesive film are cut and the adhesive's protective
covering is peeled off at room temperature. The adhesive film is
applied smoothly on the duplicate titanium strips at approximately 110
°F (43 °C) to provide additional tacking. Curing of the joint is done
by first heating it up to 250°F (121 °C) at a rate of approximately 6
®F/min (3 .3*^C/min) . At the temperature of 250 °F (121 ®C) a pressure of
40 psi (276 kPa) is applied for 60 minutes. The bonded specimen is
then cooled and removed.
The adhesive, FM 300, is a modified epoxy adhesive film supplied
with a "tricot-knit" carrier cloth that offers a good mixture of
structural and handling properties. It is manufactured by the American
Cyanamid Company as a supported film. Its product information
brochure reports a service range of -67°F to 300°F (-55 C to 150 C)
with excellent moisture and corrosion resistance.
The bonding procedure with FM 300 adhesive is described in detail
by its product information brochure [12] . Two duplicate one— inch
(2.54 cm) wide, 0.05 inch (0.127 cm) thick titanium alloy strips are
cleaned and dried to provide a grease-free surface for bonding.
Patterns of FM 300 adhesive film are cut and the adhesive's protective
covering is peeled off at room temperature. The adhesive film is then
applied smoothly on the duplicate titanium strips. Curing of the
joint is done by first heating it up to 350 °F (175°C) within a time
span of 30 to 60 minutes. At the temperature of 350°F (175°C) a
pressure of approximately 40 psi (276 kPa) is applied for 60 minutes.
16
Tho botuloa sjuH'imon i« thou coohni juul rmovoil.
Spocimon aiwonsions (width, length of overlap and adhorend
thloknoss) are in accordance with ASTM D1002 specif icationa. The
C
shear stress in the lap joints is assumed to bo nnilorm and is
calculated by dividing tho load applied to tho adhercnds by the
overlap area. The mechanical testing of tho FM 73, FM 300 and
Poly Imidesulfone-titanium specimens was performed at Clarkson College
of Technology by the methods described in the following paragraphs.
All test specimens were prepared at NASA-Langley Research Center.
Geometrical Measurements
Precision measurement of bondline thicknesses and bondlinc
lengths wore done with the aid of a microscope under a magnification
of 40X. Twenty-five or twenty-six measurements of the bondline
thickness and one measurement of the bondline length were taken along
both sides of the overlap area. The average bondline thicknesses were
used in calculating shear strains which were defined as bondline
deformation divided by the average bondline thickness. Adherend width
and thicknesses were measured with the aid of a Kanon Dial Caliper.
Single lap specimen dimensions (width, length of overlap and
bondline thickness) for the FM 73, Thermoplastic Polyimidesulfone, and
FM 300-titanlum specimens are given in tables two thru four,
respectively, in Appendix B. All dimensions listed are in both
English and Metric units.
A high precision, air cooled clip-on gage was used for shear
strain measurements. Two notches were milled 0.99 inches (2.51 cm)
apart on the overlap side of the single lap specimens so that the
17
clip-on gage could be attached in them (figure !)• The notches were
0.1 inch (0.254 cm) deep and 0.01 inch (0.0254 cm) wide. Stress
concentrations due to the notches in ^the adherends were estimated to
be negligible. Any adherend deformation that occurred within the
measuring range of the clip-on gage was calculated and subtracted from
the experimental data to obtain the adhesive strains. Such adherend
deformations were calculated with the use of the equation:
Adherend Deformation z [F(D-0)]/lE(AT)W] (22)
where
F z load applied to the adherends.
D z the distance between the milled notches in the adherends of
the single lap specimen, 0.99 inches (2.51 cm)
0 z overlap length
E - Young's Modulus for titanium, 15,000,000 psi (103,421 MPa)
AT z adherend thickness, 0.05 in. (0.127 cm)
W z adherend width over the overlap area.
The clip— on gage was calibrated with a high magnification
extensometer calibrator. Load and strain charts were obtained from a
three channel strip chart recorder. The output signal from the
clip— on gage was amplified through the mechanical testing machine s
internal amplifier system. All test specimens were stored at
approximately 70®F (21 ®C) temperature and 65 °/o humidity.
Testing Methods
Groups of twenty-four to twenty-eight single lap specimens bonded
with the model adhesives were provided by the NASA— Langley Research
Center. The mechanical testing of these single lap specimens were
18
SPRING ATTACHMENT TO
SPECIMEN TO SECURE CLIP-
ON GAGE
performed either on a Tinius-Olsen Universal testing machine or on a
Model 1331 Instron Servohydraulic testing machine depending on their
availability.
Room temperature constant strain rate tests were performed at
crosshead rates of >0.0, 0.1, 1, 5 and 10 in/min (>0.0, 0.254, 2.54,
12.7 and 25.4 cm/min) . High temperature constant strain rate tests
were performed at a crosshead rate of 5 in/min (12.7 cm/min). Room
and high temperature creep tests were performed with an initial
crosshead rate of 0.3 in/min (0.76 cm/min). Each adhesive was tested
in creep at least at four different stress levels above the elastic
limit shear stress (found from both room and high temperature constant
strain rate tests) and at three or four different temperature levels.
The FM 73 adhesive creep tests were performed at temperature levels of
70°F (21 °C), 130°F (54°C) and 180^F (82 °C). The Thermoplastic
Polyimidesulfone adhesive creep tests were performed at temperature
levels of 70°F (21 °C), 250 °F (121 °C), 350°F (177°C) and 450 F (232
°C). The FM 300 adhesive creep tests were performed at temperature
levels of 70°F (21 °C), 180 °F (82°C), 250°F (121 °C) and 300°F (148°C).
The maximum test temperature for each adhesive was determined on the
basis of product information available in the literature [10,11,121.
The high temperature creep and constant strain rate tests were
performed with the aid of an environmental chamber which was attached
to the testing machine. Deformation and load signal outputs from the
Tinius-Olsen Universal testing machine were recorded on a strip chart
recorder and the testing machine' s load chart, respectively. Similar
outputs from the Instron Servohydraulic testing machine's amplifier
were channelled to a PDP-11 Digital computer for processing of the data.
20
CHAPTER 4
RESULTS AND DISCUSSION
This chapter compares the analytical methods reviewed in Chapter
2 with the experimental constant strain ratei creep and creep-rupture
behavior of FM 73, FM 300 and Thermoplastic Poly imidesulf one
adhesives. The suitability of the proposed theoretical method for
prediction of adhesive material behavior is, therefore, assessed.
Results on FM 73 will be discussed first.
Results on FM 73 Adhesive
The constant strain rate shear stress-strain behavior of FM 73
adhesive at two different shear strain rates is shown in figure 2. A
region of linear elastic behavior followed by a region of viscoelastic
behavior is observed, thereby, suggesting that the modified Bingham
model may be used to describe the constant strain rate shear
stress-strain behavior of FM 73 adhesive. Comparison of experiment
and theory reveals that the modified Bingham model provides a good fit
O
to the experimental data* The constants (namely 0* ii> Y G) used
in fitting the model to the experimental data are also shown in figure
2 .
The elastic shear modulus, G, for FM 73 adhesive was taken to be
equal to the initial slope of the constant strain rate shear
stress-strain curve (figure 2). The value for G, as given in figure
2, is about ten times less than the value obtained by using the
relation
G zE/(2 + 2v) ^23)
21
SHEAR STRESS
vhoro
v~poisson*s ratio and,
E relaatic modulus in tension [10, 14]
-s
The authors boUove that this difference is caused by:
1) Nonuniform shear stress distribution and the presence of
tearing stresses in single lap geometry;
2) The effect of specimen geometry changes, during loading, on
adhesive deformation measurement (i*e. bending of the overlap
area) ;
3) Itivoluntary inclusion of some adherend deformation during
adhesive shear deformation measurements;
4) Inappropriate comparison of bvilk tensile measurements with
bonded (thin layer) shear measurements for adhesives with carrier
cloths.
It should bo noted that the observed differences between the
measured and calculated values is not of major concern for the
present study, since the measured values are used only in a
comparative manner to assess the effects of rate, time and temperature
on the adhesive behavior. Furthermore, the authors believe that the
measured mechanical behavior should have practical value as it
represents the actual in-service behavior of a very common joint
geometry .
Results from the constant strain rate tests for FM 73 adhesive
(.figure 2) revealed that the ultimate shear stresses are rate
dependent. This rate dependent variation in the ultimate shear
stresses with the initial elastic shear strain rate is shown in figure 3.
(
A 12% increase in the joint strength over a 500 times increase in the
23
ULTIMATE SHEAR STRESS (ksi)
A.O
3.8
3.6
3.4
shear strain rate is observed. Figure 3 also reveals that Ludwik's
equation (equation 1) provides a good fit to the experimental data.
Figure 4 shows variation of maximuifl shear strain with initial
elastic shear strain rate. Experimental data indicates a decrease in
the maximum shear strain values with an increase in the shear strain
rate as would be expected for materials that exhibit viscoelastic
behavior. An empirical equation representing the data is also shown
in figure 4.
Variation of elastic limit shear stress with initial elastic
shear strain rate is shown in figure 5. Apparently, Ludwik's equation
provides a good fit to the experimental data.
The constant strain rate shear stress 7 strain behavior of FM 73
adhesive at elevated temperatures is shown in figure 6. As expected,
the ultimate shear stress decreases with an increase in the
environmental temperature. In comparison to the room temperature
condition, the ultimate shear stress is ^ 31% , 53% and 87% lower at
130®F (54 °C), 180°F (82 °C) and 250°F (121 °C) conditions,
respectively. It can also be obseirved that the maximum shear strains
increase at elevated temperatures.
Room temperature creep behavior of FM 73 adhesive is shown in
figure 7. As may be seen, significant viscoelastic-plastic effects
(in the tertiary region) are evident at the higher stress levels.
The creep response of FM 73 at 130^F (54^C) and 180 ^F (82^C)
conditions is shown in figures 8 and 9, respectively. Higher
magnitudes of shear strains are reached at elevated temperatures,
especially at lower stress levels. The sudden plastic flow to failure
(The tertiary region) observed at room temperature (figure 7) is not
25
MAXIliUM SHEAR STRAIN (%)
10 "
INITIAL ELASTIC STRAIN RATE (%/sec)
Figure 4. Variation of Maximum Shear Strain with Initial Elastic Strain Rate
for FM 73 Adhesive.
ELASTIC LIMIT SHEAR STRESS (ksi)
30
28
26
24
22
20
18
16
14
12
10
8
6
4
Figure 5. Variation of Elastic Limit Shear Stress with Initial Elastic Strain Rate and Comparison
with Ludwik's Equation for FM 73 Adhesive.
ELASTIC LIMIT SHEAR STRESS (MPa)
SHEAR STRESS (ksi)
20 40 60 80 100
I
SHEAR STRAIN (%)
Figure 6. The Effects of Temperature on the Constant Strain Rate Stress-Strain
Behavior of FM 73 Adhesive.
SHEAR STRESS (MPa)
SHEAR STRAIN (%)
Figure 7. Room Temperature Creep Response of FM 73 Adhesive.
SHEAR STRAIN (%)
Figure 8. Creep Behavior of FM 73 Adhesive at 130°F (54°C) Condition
SHEAR STRAIN (%)
20 40 60 80
TIME (rain)
Figure 9. Creep Behavior of FM 73 Adhesive at 180®F (82®C) Condition.
evident at elevated temperatures.
Creep-rupture data for FM 73 is shown in figures 10 and 11.
Figure 10 illustrates the comparisons of Zhurkov's (modified) equation
(equation 7) with the experimental data. As may be observed, data and
theory show good agreement at elevated temperatures. The room
temperature data, however, does not show agreement with theory. The
constants and y used to fit Zhurkov's equation to the creep-rupture
data are also shown in figure 10.
Comparison of the creep-rupture data with Crochet's equation
(equation 8) is shown in figure 11. It can be seen that the behavior
is a decaying exponential one with respect to time. The plots reach
an asymptotic stress level, below which creep failures are not
expected to occur. The highest points in figure 11 (shown as solid
black figures) represent ultimate shear stress values obtained from
the highest rate shear stress-strain curves.
Figure 11 reveals that the delayed failure behavior of FM 73 can
be predicted accurately by employing either the Maxwell or modified
Bingham models. The solid and broken lines in figure 11 represent
Crochet's relations (equations 20 and 10) based on the modified
Bingham and Maxwell models, respectively. The elastic limit stresses,
(6), in equation 20 are determined by extrapolation. Since only one
"6" value corresponding to high rate testing is available at each
elevated temperature (figure 6).
The strain rates corresponding to the initial (monotonic) loading
portion of the elevated temperature creep tests are lower than those
used for instant rupture conditions of figure 11. Such low initial
strain rates are applied in the creep testing of bonded specimens to
32
Ln RUPTURE TIME (sec)
SHEAR STRESS (MPa)
Figure 10. Creep Rupture Data and Comparison with Zhurkov’s Equation for FM73
Adhesive .
SHEAR STRESS (ksi)
o □ o EXPERIMENTAL , DATA
INSTANT RUPTURE DATA FROM
CONSTANT STRAIN RATE TESTS
x(t)=A+B exp [-K(t ( t) -0 ) t]
x(t)=A+B exp [-C (x (t) ) t]
— 26
22
A (psi)
B (psi)
6 (psi)
K(psi-min)
C(psi-min) ^
T(»f)
(MPa)
(MPa)
(MPa)
(MPa-min) ^
(MPa-min)
(‘’O
3097
826'
2999
276x10'^
31x10"^
o
o
(21.4)
(5.7)
(20.7)
(1.9)x10-6
(0.2)xl0“^
(21)
1750
960
1727
120x10“^
16x10 ^
130 □
(12.1)
(6.6)
(11.9)
(0.8)xl0“^
(O.l)xlO"^
(54)
982
768
975
653x10-6
93x10-6
180 O
\^8)
(5.3)
(6.7)
(A.5)x10"6
(0.6)x10-6
(82)
120 160
TIME (min)
Figure 11. Creep Rupture Data and Comparison with Crochet's
Equation for FM 73 Adhesive.
34
SHEAR STRESS (MPa)
avoid a load overshoot. Consequently, due to the limitation in the
number of specimens available, an extrapolation procedure is applied
by assuming parallel shifts of Ludwik's equation at elevated
temperatures (figure 5) in order to obtain the necessary "0'' values.
Figure 12 shows the variation of the maximum -or safe- level of
creep stress values (below which creep failures are not expected to
occur) with environmental temperature. The maximum creep stress
values represent the asymptotic stress levels (i.e. material constants
A in figure 11) obtained with the application of Crochet's equation.
Apparently, the relation between the asymptotic shear stress level and
temperature is a linear one for the range of data shown.
Results on FM 300 Adhesive
All experimental data presented so far has been for FM 73.
Similar tests were conducted for FM 300 and similar results were
obtained. The constant strain rate shear stress-strain behavior of FM
300 adhesive at four different shear strain rates is shown in figure
13. The presence of linear elastic behavior followed by a region of
viscoelastic behavior which is terminated by plastic flow (especially
at the lower strain rate) suggests that the modified Bingham model can
be used to describe the constant strain rate shear stress-strain
behavior of FM 300 adhesive. Comparison of theory (modified Bingham
model) and experiment shows good agreement (figure 13). The constants
used in fitting the modified Bingham model to the experimental data
are also shown in figure 13.
The variation of the ultimate shear stress with initial elastic
shear strain rate for FM 300 adhesive is shown in figure 14.
35
MAXIMUM CREEP STRESS
MAXIMUM CREEP STRESS (MPa)
Figure 13. Constant Strain Rate Stress-Strain Behavior
of FM 300 and Comparison with Theory.
37
SHEAR STRESS (MPa)
ULTIMATE SHEAR STRESS (ksi)
Figure 14. Variation of Ultimate Shear Stress with Initial Elastic Strain Rate and Comparison
with Ludwik’s Equation for FM 300 Adhesive.
ULTIMATE SHEAR STRESS (MPa)
Apparently, a 2800 times increase in the shear strain rate (from
0.0751 to 213% /sec) results in -13% increase in the ultimate shear
stress. Figure 14 also reveals that liUdwik's equation (equation 1)
provides a good fit to the experimental data.
Figure 15 shows variation of maximum shear strain with the
initial elastic shear strain rate. Less scatter is observed with the
maximum shear strain values in comparison to the maximum shear stress
values. The data also indicates a decrease in the maximum shear
strain values with an increase in the shear strain rate as would be
expected for materials that exhibit viscoelastic behavior. An
empirical equation representing the data is also shown in figure 15.
Variation of elastic limit shear stress with initial elastic
shear strain rate is shown in figure 16. Apparently, Ludwik's
equation provides a good fit to the experimental data.
The effects of temperature on the constant strain rate shear
stress-strain behavior of FM 300 is shown in figure 17. In comparison
to the room temperature condition, the ultimate shear stress
is —0.5 %, 17 %,and 34 % lower at 180°F (82°C), 250°F (121°C) and 300
F (149°C) conditions, respectively. It can also be observed that,
with the exception of the 300°F (149°C) condition, the maximum shear
strains increase at elevated temperatures.
Figure 18 shows the room temperature creep behavior of FM 300
adhesive. The sudden plastic flow to failure observed in the room
temperature creep behavior of FM 73 adhesive is not present in figure
18. Creep responses of FM 300 at 180 F (82 C) , 250 F (121 C) , and 300
°F (149^C) conditions are shown in figures 19, 20 and 21,
respectively. In comparison to the room temperature behavior, higher
39
MAXIMUM SHEAR STRAIN (%)
O EXPERIMENTAL DATA
Y(%) = 37. A7 - 4.05 log (y/y')
80 — y' = 1 (%/sec)
h 2078-3
O
INITIAL ELASTIC STRAIN RATE (%/sec)
Figure 15. Variation of Maximum Shear Strain with Initial Elastic Strain Rate for FM 300
Adhesive. /
ELASTIC LIMIT SHEAR STRESS (ksi)
30
28
26
24
22
20
18
16
14
12
10
8
6
4
Figure 16. Variation of Elastic Limit Shear Stress with Initial Elastic Strain Rate and Comparison
with Ludwik's Equation for FM 300 Adhesive.
ELASTIC LIMIT SHEAR STRESS (MPa)
SHEAR STRESS (ksi)
SHEAR STRAIN (%)
Figure 17. The Effects of Temperature on the Consent
Strain Rate Stress-Strain Behavior of FM 300
Adhesive .
2
SHEAR STRESS (MPa)
SHEAR STRAIN (%)
SHEAR STRAIN (%)
TIME (min)
Figure 19. Creep Behavior of FM 300 Adhesive at 180®F (82®C) Condition
SHEAR STRAIN (%)
0 20 40 60
TIME (min)
Figure 20. Creep Behavior of FM 300 Adhesive at 250®F (12l“C) Condition
SHEAR STRAIN (%)
Figure 21. Creep Behavior of FM 300 Adhesive at 300**F (149®C) Condition.
levels of shear straios are reached at comparatively lower levels of
shear stresses. Sudden plastic flow to failure (which was not
observed with room temperature creep) is evident at elevated
temperatures.
Figure 22 shows the comparison of experimental creep-rupture data
with Zburkov's relation (equation 7). Data and theory show poor
agreement. Comparison of the creep-rupture data with Crochet's
equation based on the modified Bingham and Maxwell models (equations
20 and 10^ is shown in figure 23 along with the appropriate constants
used to fit the data. The use of the modified Bingham model results
in a slightly better fit. It should be noted that the elastic limit
stress (6) values which appear In equation (20) are extrapolated as
explained previously for FM 73 adhesive (page 32). The extrapolated
data Is shown in figure 16.
The authors believe that Crochet's equation provides a better fit
to the experimental creep-rupture data in comparison to Zhurkov's
equation. Figure 24 shows variation of the maximum - or safe - level
of creep stress values (i.e. material constants A in figure 23) with
environmental temperature. Apparently, the relation between the
asymptotic stress level and temperature is a non-linear one, for the
range shown.
Results on Thermoplastic Polyimidesulfone Adhesive
The constant strain rate shear stress-strain behavior of
Thermoplastic Polyimidesulfone at two different strain rates is shotm
in figure 25. The observed weak rate and time dependent behavior
suggests that a (non-linear) elastic model may be used to describe the
47
Ln RUPTURE TIME (sec)
SHEAR STRESS (MPa)
SHEAR STRESS (ksi)
Figure 22. Creep Rupture Data and Comparison
with Zhurkov's Equation for FM 300 Adhesive
SHEAR STRESS (ksi)
TIME (min)
Figure 23. Creep Rupture Data and Comparison with Crochet s
Equation for FM 300 Adhesive.
49
SHEAR STRESS (MPa)
MAXIMUM CREEP STRESS (ksi)
ENVIRONMENTAL TEMPERATURE (°C)
Figure 24. Variation of Maximum (Safe) Creep Stress with
Environmental Temperature for FM 300 Adhesive.
50
MAXIMUM CREEP STRESS (MPa)
SHEAR STRESS (ksi)
SHEAR STRAIN (%)\
Figure 25. Constant Strain Rate Stress-Strain
Behavior of Thermoplastic Polylmide
sulfone and Comparison with Theory.
51
SHEAR STRESS (MPa)
mechanical behavior of the adhesive* The power law models with the
use of material constants shown , provides a good fit to the
experimental data. .
Figures 26 and 27, respectively, show the variation of the
ultimate shear stresses and maximum shear strains with the initial
elastic shear strain rate* Only an 8 % increase in the joint strength
over a 100 times increase in the shear strain rate is observed
(figure 26). The maximum shear strain data indicates a relatively
constant value of maximum shear strains over a wide range of strain
rates (figure 27). Figure 26 also shows Ludwik's equation fitted to
the experimental data.
The constant strain rate shear stress— strain behavior of
Thermoplastic Polyimidesulfone at elevated t^peratures is shown in
figure 28. In comparison to the room temperature condition, the
ultimate shear stress is ^22% lower at both the 250^ (121 C) and
350 °F (177 °C) conditions; and ^44% lower at the 450°F (232°C)
condition. It can also be observed that, excluding the 350 **F (177°C)
condition, the maximum shear strains decrease at elevated
temperatures. Such a reduction in the levels of maximum shear strain
can be attributed to a change in the failure mechanism (possibly from
adhesive matrix to adhesive-fiber and/or adhesive-adherend
interfacial) of the lap joint.
Room temperature creep results for Thermoplastic Polyimidesulfone
are shown in figure 29. As can be observed, the adhesive's resistance
to creep, especially in the primary and secondary creep regions, is
characteristic of linear thermoplastics. Secondary creep rates are
much lower than those for adhesives with comparable strength values.
52
ULTIMATE SHEAR STRESS (ksl)
30
29
28
27
26
25
Figure 26. Variation of Ultimate Shear Stress with Initial Elastic Strain
Rate and Comparison with Ludwik' s Equation for Thermoplastic
Polyimidesulfone Adhesive.
ULTIMATE SHEAR STRESS (MPa)
MAXIMUM SHEAR STRAIN (%)
Figure 27. Variation of Maximum Shear Strain with Initial. Elastic Strain
Rate for Thermoplastic Polyimidesulfone Adhesive.
SHEAR STRESS (ksi)
SPECIMEN
FAILURE
20 29 T- 3
2032T-4
(%/sec)
(°F)
(”C)
24.6
70
21
o
59
250
121
A
86
350
177
□
47
232
O
SHEAR STRAIN (%)
The Effects of Temperature on the
Constant Strain Rate Stress-Strain
Behavior of Thermoplastic Polyimide
sul f one Adhe s ive .
SHEAR STRESS (MPa)
Figure 29. Room Temperature Creep Response of Thermoplastic Polyimidesulfone
Adhesive.
The presence of a tertiary creep region is evident. It should also be
noted that the level of initial shear strain (y^,) reached for a given
level of shear stress appears to control the oncoming creep process
(compare creep behavior at 3273 psi - 22.6 MPa and, 3351 psi - 23
MPa) .
The creep results of Thermoplastic Polyimidesulfone at 250 °F (121
°C), 350 °F (177°C) and 450®F (232°C) conditions are shown in figures
30, 31 and 32, respectively. A decrease in the levels of creep
strains is observed with increasing temperatures, especially at the
(232^0) condition where an almost constant level of shear
strain is evident. The reason for such a decrease in the levels of
shear strain becomes obvious when one examines the post-failure
surfaces of the specimens (figure 33). The extent of interfacial
separation, especially adhesive-adherend (and also adhesive-fiber)
separation increases with an increase in the environmental
temperature. The gradual darkening of the post-failure overlap areas
with increasing temperatures (in figure 33) is an indication of
adhesive matrix layers which have separated from the adherend and/or
the carrier cloth. Consequently, interfacial separations result in
failures without appreciable adhesive deformation.
Creep-rupture data for Thermoplastic Polyimidesulfone is shown in
figures 34 and 35. Figure 34 shows the comparison of Zhurkov's
(modified) equation (equation 7) with the experimental data. As may
be observed, data and theory show poor agreement. The discrepancy
between data and theory is greater at the room temperature and 450°F
(232°C) conditions. The constants Uq and Y used to fit Zhurkov s
equation to the creep-rupture data are also shown in figure 34.
57
SHEAR STRAIN (%)
40
35
30
25
20
15
10
SHEAR STRAIN (%)
SPECIMEN FAILURE
T = 2542 psi
° = 17.5 MPa
2045T-4
350*F
177°C
T - 1905 psi (13.1 MPa) SPECIMEN REACHED lOlSTRAIN AFTER 50.6 HRS. 2096T-2
t° - 2274 psl (15.7 MPa) SPECIMEN FAILED AFTER 15.5 HRS. AT 7.5% STRAIN 2046T 1
TIME (min)
Figure 31. Creep Behavior
Condition.
of Thermoplastic Poly imidesulf one Adhesive at 350*F (177*0
SHEAR STRAIN (%)
18.77
SPECIMEN FAILURE
T = 985 psi
° = 6.79 MPa
20A6T-4
T = 450°F
= 232'’C
= 375 psl (2.58 MPa) SPECIMEN REACHED 17.9% STRAIN AFTER 36 HRS. 20A7T-1
17.86
Figure 33. Typical Fracture Surfaces of Tltanium-Polylmlclesulfone Specimens Exhibiting
the Effects of Increasing Temperature.
|i|
1
m
Ln RUPTURE TIME (sec)
SHEAR STRESS (MPa)
Figure 34. Creep Rupture Date and Compariaon with Zhurkov’a Equation for Thermo
plastic Polyimidesulfone Adhesive.
Figure 35 shows the comparison of creep-rupture data with
Crochet's equation (equation lo) based on the Maxwell model.
Apparently, Crochet's equation provides oa better fit to the
experimental data in comparison to Zhurkov's equation. The appropriate
constants used in fitting Crochet's equation (based on the Maxwell
model) to the creep-rupture data are also shown in figure 35.
Figure 36 shows variation of the maximum -or safe- level of creep
stress values(below which creep failures are not expected to occur)
with environmental temperature. The maximum creep stress values
represent the asymptotic shear stress levels (i.e. material constants
A in figure 35) obtained with the application of Crochet's equation.
A sharp decrease in the level of maximum creep stress is observed for
temperatures above 350^ (177®C).
A comparison of the mechanical properties of FM 73 , FM 300 and
Thermoplastic Polyimidesulfone adhesives at comparable testing
conditions is shown in Table 1 (page 66) . As may be observed from
Table 1, the epoxy adhesives, FM 73 and FM 300, are clearly inferior
to the polyimide adhesive. Thermoplastic Polyimidesulfone, in high
temperature strength retention. The creep resistance of the polyimide
adhesive is higher than that for the epoxy adhesives, especially at
elevated temperatures.
63
SHEAR STRESS (ksi)
Figure 35. Creep Rupture Data and Comparison with Crochet's
Equation for Thermoplastic Polyimidesulf one
Adhesive.
64
SHEAR STRESS (MPa)
MAXIMUM CREEP STRESS (ksi)
ENVIROMMENTAT. TEMPERATURE (°C)
ENVIRONMENTAL TEMPERATURE (°F)
Variation of Maxinuim (s'afe) Creep Stress with
Environmental Temperature for Thermoplastic
Polyimidesulfone Adhesive.
MAXIMUM CREEP STRESS (MPa)
Property and
Test Condition
Elastic Shear
Rodulus, G
T=70°F (21“C)
Maximum Shear
Stress and Strain
Under Constant
Strain Rate
Conditions at
70°F (21°C)
Maximum Shear
Stress and Strain
Under Constant
Strain Rate
Conditions at
250°F (121°C)
Room Temperature
Creep Strain
at Comparable
Creep Stresses
Maximum Creep
Strain at
180°F (82°C)
Maximum Creep
Strain at
250 F (1210
Maximum (Safe)
Level of Creep
Stress at 70 Pf
(21°C)
FM 73
Adhesive
,12.7 ksi
(87 .6 MPa)
3,469 psi
(23.9 MPa)
58.5 %
500 psi
(3.45 MPa)
107 %
85.25 %
3,499 psi
(24.1 MPa)
130 %
1,104 psi
(7.6 MPa)
3,097 psi
(21.4 MPa)
FM 300
Adhesive
19.7 ksi
(436.1 MPa)
4,300 psi
(29.7 MPa)
31.2 %
(0.75 %/sec) (1.94 %/sec)
3,652 psi
(25.2 MPa)
72.8 %
(2.36 %/sec) (384 %/sec)
32.01%
3,535 psi
(24.4 MPa)
65 %
2,925 psi
(20.2 MPa)
125 %
2,200 psi
(15.2 MPa)
3,675 psi
(25.3 MPa)
Thermoplastic
Poly imidesul f one
25.1 ksi
(173.0 MPa)
3,947 psi
(27.2 MPa)
59.6 %
( 1 .35 %/ sec)
3,450 psi
(23.8 MPa)
37.5 %
(59.1 % / sec)
56.76 %
3,423 psi
(23.6 MPa)
37.76 %
3,117 psi
(21.5 MPa)
3,267 psi
(22.5 MPa)
Maximum (Safe)
Level of Creep
Stress at
180°F (82 C)
Maximum (Safe)
Level of Creep
Stress at
250 F (12rC)
982 psi
(6.8 MPa)
( )
2,800 psi
(19.3 MPa)
2,07 5 psi
(14.3 MPa)
( )
2,765 psi
(19.1 MPa)
Table 1. Mechanical Properties of FM 73, FM 300 and Thermoplastic
Polyimidesulfone Adhesives at Comparable Testing Conditions.
66
V. CHAPTER 5
CONCLUSIONS
0
The present investigation was concerned with the development of a
practical method for characterizing structural adhesives in the bonded
lap shear mode. The validity of this proposed characterization method
was evaluated with the use of experimental constant strain rate, creep
and creep-rupture data. Three different model adhesives were used in
the bonded lap shear mode for this purpose. Elevated temperature
behavior was also studied#
Based on the experimental data from the three adhesives studied,
the following conclusions can be madet
1) It is possible to describe the constant strain rate
shear stress— strain behavior of structural adhesives
in the bonded form by using viscoelastic or nonlinear
elastic relations.
2) Ludwik's equation provides an adequate description of
the rate dependent ultimate and elastic limit shear stresses.
3) Crochet’s equation provides a better description of
the temperature dependent creep-rupture data in
comparison to Zhurkov’s equation.
4) The epoxy adhesives, FM 73 and FM 300, are clearly inferior
to the polyimide adhesive. Thermoplastic Polyimidesulfone,
in high temperature strength retention.
The authors’ suggestions for future work include correlation of
bulk adhesive properties to single lap shear data. The effects of
temperature on the rate dependent behavior and viscosity of adhesives
67
should also be studied as a continuation of the present
investigation. Results from such a study can be correlated to those
from the constant strain rate tests conducted at elevated temperatures
herein. Future efforts in the study of the effects of polymer
molecular weight on adhesive mechanical behavior should also prove to
be worthwhile.
68
REFERENCES
1. Brinson, H.F. , "The Viscoelastic-Plastic Characterization of a
Ductile Polymer," Deformation and Fracture of High Polymers, H.
Kaush et. al., eds» , Plenum Press, NY, 1974«
2. Brinson, H.F. , Renieri, M.P. , and Herakovich, C.T., "Rate and Time
Dependent Failure of Structural Adhesives," Fracture Mechanics of
Composites, ASTM STP 593, 1975, pp. 177-199.
3. Sancaktar, E. and Brinson, H.F. , "The Viscoelastic Shear Behavior
of a Structural Adhesive,” Adhesion and Adsorption of Polymers,
12-A, Plenum Publishing Corp., 1980.
4. Chase, K.W. and Goldsmith, W. , "Mechanical and Optical
Characterization of an Anelastic Polymer at Large Strain Rates and
Strain," Experimental Mechanics, p. 10, Jan. 1974.
5. Sancaktar, E. and Padgilwar, S., "The Effects of Inherent Flaws on
the Time and Rate Dependent Failure of Adhesively Bonded -Joints,
Transactions of the ASME Journal of Mechanical Design, Vol. 104,
No. 3, pp. 643-650, (1982).
6. Thorkildsen, R.L. , Engineering Design for Plastics, E. Baer (Ed.),
Reinholt Book Co., New York (1964) , p. 295.
7. Zhurkov, S.N. , "Kinetic Concept of the Strength of Solids,"
Internl. J. of Frac. Mech., 1 (4), 331-323 (1965).
8. Crochet, M. J. , "Symmetric Deformations of Viscoelastic-Plastic
Cylinders," J. of Appl. Mech., 33, p.326, 1966.
9. Bugel, T.E., Norwalk, S. and Snedeker, R.H., "Phenoxy Resin - A New
Thermoplastic Adhesive," Adhesion, Ed. Eley , D.D. , pp. 87-94,
Oxford University Press, 1961.
10. St. Clair, T.L. and Yamaki, D.A. , "A Thermoplastic Polyimide-
sulfone," NASA Langley Research Center, Technical Memorandum
No. 84574, Nov. 1982.
11. FM 73 Adhesive Film Product Information Brochure. American
Cyanamid Company, Bloomingdale Products, Havre de Grace, Maryland
21078.
12. FM 300 Adhesive Film Product Information Brochure. American
Cyanamid Company, Bloomingdale Products, Havre de Grace, Maryland
21078.
13. Goland, M. and Reissner, E. , "The Stresses in Cemented Joints," J.
Applied Mechanics, Vol. II, No. 1, March 1944.
69
14. Mall, S., Johnson, W.S. and Everett, R.A., Jr., "Cyclic Debonding
of Adhesively Bonded Composites," NASA Langley Research Center,
Technical Memorandum No. 84577, Novo. 1982.
15. Nadai, A. and Manjoine, M.J., Journal of Applied Mechanics, vol. 8
p. A82, 1941.
16. Slonimskii, G.L. , Askadskii, A.A. and Ka 25 antseva, V.V. , "Mechanism
of Rupture of Solid Polymers," Polymer Mech., 5 (19771.
17. Brinson, H.F., Griffith, W.I. and Morris, D.H. , "Creep Rupture of
Polymer-Matrix Composites," Experimental Mechanics, Sept. 1981, pp.
329-335.
18. Reiner, M. , Advanced Rheology, H.K. Lewis and Co. Ltd., London,
208, 1971.
19. Schapery, R.A. , "On the Characterization of Nonlinear Viscoelastic
Materials," Poly. Eng. and Sci., Vol. 9, No.. 4, July 1969, pp.
295-310.
20. Cartner, J.S. and Brinson, H.F. , "The Nonlinear Viscoelastic
Behavior of Adhesives and Chopped Fiber Composites, Virginia
Polytechnic Institute Report No. VPI-E-78-21, August 1978.
21. Greenwood, L. , "The Strength of a Lap Joint," Aspects of Adhesion,
Vol. 5, 1969, pp. 40-50.
70
APPENDIX A
CALCULATION OF OVERLAP EDGE
STRESS CONCENTRATION FACTORS.
71
The quantity reported as the "lap shear strength,” is the
breaking load per unit bond area* However, failure initiation in
materials is a localized phenomenon that is dependent on the ma x i m um
stresses at a point reaching some critical value. The standard lap
shear specimen shows a variation in the shear stress within the
adhesive layer and, in fact, can exhibit a singular behavior at the
bond termination.
Volkersen [21] analyzed the stress distribution in single lap
joint geometry under a load due to stretching of the adherends while
ignoring the tearing stresses at the free ends* He considered the
case where two identical elastic adherends of uniform thickness, t and
a Young's moduli of E, bonded over a length of 2C by an adhesive of
thickness m which was assumed to behave like an elastic solid with
shear modulus Gq^* The bonded faces of the adherends were assumed to
be parallel so that the thickness m was considered constant*
Volkersen' s analysis showed that when the bonded members are in pure
tension, the shear stress in the adhesive layer is a maximum at each
end of the overlap* He compared the maximum shearing stress at the
end of the overlap with the mean stress to evaluate the stress
concentration factor.
n^z(/ coth (oO
(24)
where
a
( 25 )
72
Goland and Reissner [13] , in their analysis of the single lap
joint geometry, showed the existence of high stress concentrations
near each end of the overlap region, in the form of peak normal
0
(tearing) and shear adhesive stresses. The magnitudes of these
stresses depend on
1) the flexibility of the adhesive,
2) the length of overlap,
3) Young's moduli of the adherends,
4) the thickness of the adherends, and
5) the thickness of the adhesive layer*
Goland and Reissner categorized adhesively bonded joints into two
approximate groups to simplify their analysis.
In the first approximation, the work of the normal (or'^,) and shear
( adhesive stresses were neglected in comparison to the work of the
adherend stresses (of, andXwv/) which were assumed to be continuous
y
across the adhesive layer. Adhesive layers were considered inflexible
for joints satisfying the conditions
(jn/E ) « (t/E), Cni/G ) « (t/G) . (26)
In the second approximation, the work of the adherend stresses was
ignored in comparison to the work of adhesive stresses andr^, for
joints satisfying the conditions
(m/E ) » (t/E) or (m/G ) » (t/G). (27)
Such joints were considered to have flexible adhesive layers.
73
The single lap joints used for the present investigation can be
considered to have flexible adhesive layers. Based on the Goland and
Reissner analysis [13]} the stress concentration factor^
flexible joints can be calculated as
n = (l/4)l(BC/tHl+3k)coth(BC/t)+3(l-k)j
GR
( 2 »)
where
and
also:
(29)
(30)
(31)
(32)
B =\KRG t/Em) ,
^ a
k= (cosh(UC))/lco8h(UC)+2\^sinh(UC)) ,
UC =\/l3(lV')/2j (C/t)lF/(tWE)jK^
y = [6E^t/Bnf-/-i*
a
v= poisson's ratio for the adherends (0.30)
F = load applied to the adherends
E = elastic modulus for the adhesive
a
ra = adhesive layer thickness
W = width of the adherend.
It should be noted that this approach again neglects the presence of
tearing stresses in the adhesive layer. A more accurate analysis
should include the effect of the tearing stresses as described below.
Since both normal (0^*,)^^^^^ .shear stresses exist at the
overlap edge, principal (normal), and maximum shear, stresses
need to be calculated for the assessment of the failure condition.
These stresses can be calculated with the use of equations
74
( 33 )
= (o' ) /2
©max'
[((o' ) /2)'
© max'
and.
^max
[((ol)^ /2)‘
o max
((T )
o max
(34)
where (T ) is given by equation (28) and (o'!) is determined by
Oina©c « max
the method of Goland and Reissner [13] as
(O = (F/tW)(C/t)^ [(L^k/2)(8inh(2L)-sin(2L))/(8inh(2L)+8in(2L))
° meix
+Lk • ( CO 8h( 2L) +C0 s( 2L) ) / ( 8inh( 2L) + 8in( 2L) ) J (3b)
where
L = (YC/t),
k' = (V^jCW/Ft),
(36)
(37)
and,
Vq = (kF/W)l3F(l-v^)/(tEW)jV^ (38)
More accurate stress concentration factors n^ and n^^can now be
calculated by dividing and respectively by the average shear
stress (T r F/A) .
Application of the four methods discussed above for calculating
the stress concentration factor, n, for single lap joints bonded with
FM 73, FM 300 and Thermoplastic Polyimldesulfone adhesives under
constant strain rate loading yields the following results;
75
Adhesive
Specimen
“V
n
GR
V
"o'
FM 73
SJ-9-4
1.8828
2.4552
2.7458
3.9752
SJ-10-2
2.1493
2.8000
3.1404
4.5404
FM 300
2078-3
2.0283
2.6549
2.9721
4.3081
2083-3
2.0017
2.6264
2.9404
4.2626
2078-2
1.9937
2.6126
2.9242
4.2376
2083-4
1.8189
2.4305
2.7164
3.9295
Thermoplastic
2027T-1
2.3917
3.0312
3 .4023
4.9476
Poly imidesulf one 2029T-3
2.3684
3 .0296
3.3992
4.9408
where for
FM 73 adhesive
E^= 337,068.87 psi [14]
V =0.40 [14] ,
FM 300 adhesive
E, = 337,068.87 psi [14]
v= 0.40 [14]
and,
Ther moplastic Poly imidesulf one
E = 719,000 psi [10]
3
V =0.38 [10].
76
APPENDIX B
TABLES OF SINGLE LAP SPECIMEN DIMENSIONS
77
TABLE B-2
Dimensions for Single Lap Specimens Bonded
FM 73 Adhesive.
with
Specimen ]
No.
Jondline Thickness
Inches (CM)
Overlap Length
Inches (CM)
Overlap Width
Inches (CM)
SJ-5-1
0.00540(0.01372)
0.51969(1.32001)
0.99402(2.52481)
SJ-5-2
0.00470(0.01194)
0.52362(1.32999)
0.99410(2.52501)
SJ-5-3
0,00480(0.01219)
0.51969(1.32001)
0.99420(2.52527)
SJ-5-4
0.00520(0.01321)
0.53150(1.35001)
0.99475(2.52667)
SJ-6-2
0.00550(0,01397)
0.50886(1.29250)
0.99783(2.53449)
SJ-6-4
0.00740(0.01880)
0.50689(1.28750)
0.99616(2.53025)
SJ-7-1
0.00460(0.01168)
0.51083(1.29751)
0.99700(2.53238)
SJ-7-2
0.00500(0.01270)
0.51280(1.30251)
0.99050(2.51587)
SJ-7-3
0.00480(0.01219)
0.51870(1.31750)
0.98885(2.51168)
SJ-7-4
0.00480(0.01219)
0.52264(1.32751)
0.98850(2.51079)
SJ-8-1
0.00519(0.01318)
0.51969(1.32001)
0.99140(2.51816)
SJ-8-3
0.00501(0.01273)
0.51772(1.31501)
0.99265(2.52133)
SJ-9-2
0.00500(0.01270)
0.51969(1.32001)
0.99409(2.52499)
SJ-9-3
0.00550(0.01397)
0.51870(1.31750)
0.99499(2.52600)
SJ-9-4
0.00650(0.01651)
0.51673(1.31249)
0.99595(2.52971)
SJ-10-1
0.00540(0.01372)
0.53150(1.35001)
0.99675(2.53175)
SJ-10-2
0.00480(0.01219)
0.52559(1.33500)
0.99425(2.52540)
SJ-10-3
0.00474(0.01204)
0.52067(1.32250)
0.99270(2.52146)
SJ-10-4
0.00593(0.01506)
0.51378(1.30500)
0.99240(2.52070)
78
Dimensions for Single Lap Specimens Bonded with
Thermoplastic Polyimidesulfone Adhesive,
Specimen
No.
Bondline Thickness
Inches (CM)
Overlap Length
Inches (CM)
Overlap Width
Inches (CM)
2026T-4
0.00998(0.02535)
0.50500(1.28270)
0.99500(2.52730)
2027T-1
0.00839 (0.02131)
0.52580(1.33553)
0.98540(2.50292)
2027T-2
0.00775(0.01969)
0.52006(1.32095)
0.98700(2.50700)
2028T-4
0.00970(0.02464)
0.53800(1.36652)
0.98100(2.49174)
2029T-1
0.00897(0.02278)
0.51400(1.30556)
0.98500(2.50190)
2029T-2
0.00919(0.02334)
0.51500(1.30810)
0.98400(2.49936)
20 29 T- 3
0.00840(0.02134)
0.52099(1.32331)
0.98600(2.50444)
2030T-3
0.00889(0.02258)
0.53481(1.35842)
0.98540(2.50292)
2030T-4
0.00987(0.02507)
0.53684(1.36357)
0.98540(2.50292)
2031T-1
0.00869 (0.02207)
0.52400(1.33096)
0.99100(2.51714)
2031T-2
0.00796(0.02022)
0.52500(1.33350)
0.98900(2.51206)
2031T-3
0.00797(0.02024)
0.52900(1.34366)
0.99400(2.52476)
2032X-3
0.00909(0.02309)
0.52710(1.33883)
0.97913(2.48699)
2032T-4
0.00894(0.02271)
0.53144(1.34986)
0.97125(2.46698)
2044T-1
0.01021(0.02593)
0.52953(1.34501)
0.98450(2.50063)
2044T-2
0.00983(0.02497)
0.52953(1.34501)
0.98455(2.50076)
2044T-3
0.00952(0.02418)
0.53642(1.36251)
0.98675(2.50635)
2044T-4
0.01058(0.02687)
0.53740(1.36500)
0.99315(2.52260)
2045T-1
0.01135(0.02883)
0.507.87(1.28999)
0.99675(2.53175)
2045T-4
0.01094(0.02779)
0.50984(1.29499)
1 .00300(2.54762)
2046T-1
0.00949(0.02410)
0.48819(1.24000)
0.99078(2.51658)
2046T-2
0.00871(0.02212)
0.50394(1.28001)
0.98980(2.51409)
2046T-3
0.00870(0.02210)
0.51378(1.30500)
0.98775(2.50889)
2046T-4
0.00859(0.02182)
0.52657(1.33749)
0.98675(2.50635)
2047T-1
0.01084(0.02753)
0.54134(1.37500)
0.99150(2.51841)
2047T-2
0.00895(0.02273)
0.53150(1.35001)
0.97750(2.48285)
2047T-3
0.00931(0.02365)
0.52362(1.32999)
0.97680(2.48107)
79
TABLE B-4
Dimensions for Single
FM 300 Adhesive.
Lap Specimens Bonded with
Specimen
No.
Bondline Thickness
Inches (CM)
Overlap Length
Inches (CM)
Overlap Width
Inches (CM)
2078-2
0.00575(0.01460)
0.53346(1.35499)
0.99065(2.51625)
2078-3
0.00574(0.01458)
0.54230(1.37744)
0.98800 (2.50952)
2078-4
0.00644(0.01636)
0.54530(1.38506)
0.98800(2.50952)
2079-i
0.00678(0.01722)
0.52756(1.34000)
0.99225(2.52032)
2079-2
0.00596(0.01514)
0.52559(1.33500)
0.99075(2.51651)
2079-3
0.00647(0.01643)
0.51969(1.32001)
0.99050(2.51587)
2079-4
0.00612(0.01554)
0.51378(1.30500)
0.99150(2.51841)
2080-2
0.00629(0.01598)
0.53839(1.36751)
0.98800(2.50952)
2080-3
0.00594(0.01509)
0.53839(1.36751)
0.98650(2.50571)
2080-4
0.00657(0.01669)
t
0.53642(1.36251)
0.98575(2.50380)
2081-1
0.00524(0.01331)
0.52953(1.34501)
0.98250(2.49555)
2081-2
0.00535(0.01359)
0.53346(1.35499)
0.97925(2.48730)
2081-3
0.00503(0.01278)
0.53543(1.35999)
0.97775(2.48348)
2081-4
0.00491(0.01247)
0.53740(1.36500)
0.97825(2.48475)
2082-2
0.00592(0.01504)
0.53150(1.35001)
0.99000(2.51460)
2082-3
0.00601(0.01527)
0.52953(1.34501)
0.99000(2.51460)
2082-4
0.00588(0.01494)
0.52539(1.33449)
0.99325(2.52285)
2083-1
0.00611(0.01552)
0.54528(1.38501)
0.98975(2.51396)
2083-2
0.00516(0.01311)
0.53743(1.36507)
0.98775(2.50888)
2083-3
0.00562(0.01427)
0.52953(1.34501)
0.98375(2.49873)
2083-4
0.00647(0.01643)
0.51575(1.31001)
0.98975(2.51397)
2084-1
0.00501(0.01273)
0.54921(1.39499)
0.98475(2.50127)
2084-2
0.00588(0.01494)
0.55118(1.40000)
0.98650(2.50571)
2084-3
0.00573(0.01455)
0.55118(1 .40000)
0.98950(2.51333)
80
1. Report No.
NASA CR-172237
2. Government Accession No.
4. Title and Subtitle
material characterization of stroctoral adhesives in
THE LAP SHEAR MODE
7. Author(s)
Steven C. Schenck and Erol Sancaktar
9. Performing Organization Name and Address
riarkson College of Technology
Department of Mechanical and Industrial Engine.erxng
Potsdam, NY 13676
I 12. Sponsoring Agency Name and Address
National Aeronautics and Space Administration
Washington, DC 20546
3. Recipient's Catalog No. «
5. Report Date
September 1983
6. Performing Organization Code
8. Performing Organization Report No.
10. Work Unit No.
11. Contract or Grant No.
NAGI-284
13. Type of Report and Period Covered
Contractor Report
14, Sponsoring Agency Code
16. Abstract
A general method for fom^'orsLi-emJSLraS ^prLches
is proposed. Two approaches in t , . ^ Tndwik's and Zhurkov's equations
ar/ Jed. The see.l-emplrlcal ^J^rfaSSferin t^e c Jtant strain raje and
to describe respectxvely , th , . nf the temperature effects,
constant stress loading modes with ® adhesive shear stress-strain behavior
The theoretical approach is used ^^^^J^^'^^^^titutive equations. Three
with the use of viscoelastic or nonlinear ^^de with titanium
different model adhesives JL^developed at NASA Langley Research
adherends. These ^ for possible aerospace applications.
Center) are currently considered by NASA tor poss the generality of the
Use of different model adhesives helps in assessment of the generality
method.
1 17. Key Words (Suggested by Authof(s))
adhesive creep rupture
visco-elastic single lap specimen
nonlinear-elastic thermoplastic adhesive
rate effects high temperature
creep Bingham model
18. Distribution Statement
Unclassified - Unlimited
Subject Category 27
I 19. Security Oassif. (of this report)
Unclassified
20. Security Classif. (of this page)
Unclassified
21. No. of Pages
88
22 Ptice
AOS
N-305
For sale by the National Technical Information Service, Springfield. Virginia 22161
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